Fatigue characterization of fabricated ship details Phase 2

(1)

,4,

SSC-346

FATIGUE CHARACTERIZATION

OF FABRICATED SHIP DETAILS

(Phase 2)

This document has been approved

for public release and sale; its distribution is unlimited

SHIP STRUCTURE COMMITTEE

(2)

SHIP STRUCTURE COM1IUEE

The SHIP STRUCTURE COMMI I hE is constituted to prosecute a research program to improve the hull structures of ships and other marine structures by an extension of knowledge pertaining to design, materials, and methods of construction.

RADM J. D. Sipes, USCG, (Chairman)

Chief, Office of Marine Safety, Security and Environmental Protection

U. S. Coast Guard

Mr. Alexander Malakhofl Director, Structural Integrity

Subgroup (SEA 55Y)

Naval Sea S'stoms Command

Dr. Donald Llu

Senior Vice President American Bureau of Shipping

CONTRACTING OFFICER TECHNICAL REPRESENTATIVES Mr. William J. Siekierka

SEA 55Y3

Naval Sea Systems Command

SHIP STRUCTURE SUBCOMMITTEE

The SHIP STRUCTURE SUBCOMMITTEE acts for the Ship Structure Committee on technical matters by providing technical coordination for determinating the goals and objectives of the program and by evaluating and interpreting the results in terms of structural design, construction, and operation.

AMRlCAN;U- _

."I

Mr. Stephen G. Arntson (Chairman) Mr. John F. Conlon

Mr. William Hanzalek Mr, Philip G. Rynn

MILITARY SEALIFT COMMAND

Mr. Albert J. Attermeyer Mr. Michael W. Tourna Mr. Jeffery E. Beach MARITIME ADMINISTRATION Mr. Frederick Seibold Mr. Norman O. Hammer Mr. Chao H. Lin Dr. Walter M. Maclean

SHIP STRUCTURE SUBCOMMITTEE LIAISON MEMBERS U. S. COAST GUARD ACADEMY

LT Bruce Mustain

LI, S. MERCHANT MARINE AÇJ,DEMY

Dr. C. B. Kim

U.S. NAVAL ACADEMY Dr. Ramswar Bhattacharyya

STATE UNIVERSITY OF NEW YORK MARITIME COLLEGE

Dr. W. R. Porter

WELDING RESEARCH COUNCIL

Mr. H. T. Haller

Associate Administrator for

Ship-building and Ship Operations

Maritime Administration

Mr. Thomas W. AlIen Engineering Officer (N7)

Military Sealift Command

CDR Michael K. Parmelee, USCG, Secretary, Ship Structure Committee

U. S. Coast Guard

Mr. Greg D. Woods SEA 55Y3

Naval Sea Systems Command

NAVAL S.EA SYSIEMS COMMANQ Mr. Robert A. Sielski Mr. Charles L. Null Mr. W. Thomas Packard Mr. Allen H. Engle U. S. COAST GUARD CAPT T. E. Thompson

CAPT Donald S. Jensen

CDR Mark E. NOII

NATIONAL ACADEMY OF SCIENCES

-MARIN E BOARD

Mr. Alexander D. Stavovy

NATIONAL ACADEMY OF SCIENCES -COMMITTEE ON MARINE STRUCTURES Mr. Stanley G. Stiansen

SOCIETY OF NAVAL ARCHITECTS AND MARINE ENGINEERS

-HYDRODYNAMICS COMMITTEE Dr, William Sandberg

AMERICAN IRON AND STEEL INSTITUTE Mr. Alexander D. Wilson

(3)

Member Agencies:

United States Coast Guard Naval Sea Systems Command Maritime Administration Amercan Bureau of Shipping Military Sealift Command

Ship

Structure

Cornmittee

An Interagency Advisory Committee

Dedicated to the Improvement of Manne Structures

December 3, 1990

FATIGUE CHARACTERIZATION OF FABRICATED

SHIP DETAILS (PHASE 2)

A basic understanding of fatigue characteristics of fabricated

details

is necessary to ensure the continued reliability and

safety of

ship structures.

Phase

1

of this study (SSC-318)

provided a fatigue design procedure for selecting and evaluating

these details.

In this second phase,

an extensive series of

fatigue tests were carried on structural details using variable

loads

to

simulate

a

vessel's

service history.

This

report

contains the test results as well as fatigue predictions obtained

from available analytical models.

Address Correspondence to:

Secretary, Ship Structure Committee U.S. Coast Guard (G-Mm)

2100 Second Street S.W. Washington, D.C. 20593-0001 PH: (202) 267-0003 FAX: (202) 267-0025 SSC- 34 6 SR- 1298

J) SIPES

Rear Admiral, U.S. Coast Guard

(4)

Technical Report Documentation Page 1. R.pert No.

SSC-346

2. Gor.rnm.ne Acc.ss,en N0. 3. R.cip..nts Catalog No.

1. TitI. and Subtitl

Fatigue Characterization of Fabricated Ship Details for Design - Phase II

5. Report Dote

May 1988

6. P.rforming Organiiat,orr Cod.

Ship Structure Committee

8. P.rforrning Orgonizotion R.port No.

SR- 12 98

7. Author's)

S. K. Park and F. V. Lawrence

9. Performing Orgen.ze?,en Nom. end Address

Department of Civil Engineering

University of Illinois at Urbana-Champaign

205 N. Mathews Avenue

Urbana, IL 61801

10. Work Uni? No. (TRAIS)

II. Contract or Grant No.

DTCG 23-84-C-20018

13. lype f Report end P.riod C..r.d

Final Technical Report

12. Sponsorng Agency Nome and Ad.ss

U. S. Coast Guard 2100 2nd Street S.W.

Washington, DC 20593 14. Sponsoring Agency Cod.

G-M

IS. Suppl.n'entory Notes

The U.S.C.G. acts as the contracting office for the Ship Structure Committee.

16. Abstract

The available analytical models for predicting the fatigue behavior _of

weldinents under variable amplitude

load histories were compared using test results for weidments subjected to the SAE bracket and transmission _variable

load amplitude histories. _{Models based on detail S-N diagrams}

such as the Munse

Fatigue Design Procedure (MFDP) _{were found to perform well except when the}

history had a significant _{average mean stress.}

Models based on fatigue crack propagation alone were generally conservative, while a model based _{upon estimates} of both fatigue crack initiation

and propagation (the I-P Model) performed the best.

An extensive series of fatigue _{tests was carried out on welded}

structural

details commonly encountered in ship construction using a variable load _history which simulated the service _{history of a ship.}

The results from this study

showed that linear cumulative _{damage concepts predicted the}

test results, but the

importance of small stress _{range events was not studied because}

events smaller

than 68 MPa (10 ksi) _{stress range were deleted from the developed}

ship history to

reduce the time required for _testing.

An appreciable effect of _{mean stress was} observed but the results did not verify the existence of _{a specimen-size effect.}

Baseline constant-amplitude S-N _{diagrams were developed for five}

complex

ship details not commonly studied _{in the past.}

17. Key Words

Fatigue, Ship Structure Details, Design, Reliability, Loading History, Variable Load Histories

18. Distribution Stat.,nenl

Document is available to the U.S. public, the National Technical Information Service,

Springfield, VA 22151

19. Security Clossif. (et thu report)

Unclassified

- 20. Security Classif. (of this page)

Unclassified

21. No. of Pagea

201

22. Price

(5)

Appisilmils C.av.iskas I. Mstik Musai..

WI.. Y.. 5Mw

Multiply by

le Pld

$S1

METRIC CONVERSION FACTORS

M

a

-R MAIS (wsight) _____

E

sS 25 __________

.

5.03 hiI.,u

E

_tMs. s., iss... t -125501W a . VOLUME WSIpOs.5 I )IIlIi4«p sI ubI..ss.i il s.lIiIiis.i Ml -fluid a.c.0 21 iÑIlilil.,. M . C15 5.24 IÓIS(S I

E

i. pat. .4

lis

I -qNsiti 5.21 Ill.i. I øslIsss 3.5 lii«. I

-__ a CubiC Isst 0.03 CubiC .41511 Ui3 cubic y.id. 5.75 CubiC 5515v.

a

-TEMPERATURE (es.ct)

______

-FIkiSidiSit III (stt Culsiul

=

M tI51Çela4«S Iuh.òClIil tsispstltas

-32) II

ZI ..j.II3. I.. 011M oau

CWI.,I5J lI.C. e. 315$ hic. P,ci.

S6,

ol *.,lii - Me.i.11. Pvc. 12.2$. SO Cuijiul N... C13.IO:255.

App,.xi.ale C.avets4e.s is. Mehl. M..s.i.s

$1

WI.. V.. Ssaa Multiply by T. Had SVUISI LINSTN AIEA Cpi' 5 CpR4h.4N5 SIS 54 I.u5 IU 2 s., isvs

ii

s,ICM ys.s Is.' s hlI..IS($ 5.4 be h.cs. (IS SOS s( 2.1 c MASS (walsh s 0.s* C.. SI kltsCMa 2.2

pis

t Ms... lION kil 1.1 OLUMI 5) .IIlilit.v. 0.53 huid ss.s.s It sc I lIess LI paw p' I lIsis IN spsisi I huis 0.21 NIls.. as) Cubic 541515 21 cubic 1551 fI cubic .411«s 1.3 cubIc ylids pl3 1.3 cubIc ylids pl3

TEMPERATURE I....)TEMPERATURE I....)

Calsius S/I (di.. habisiMsil P

t«as

.

12) siMiN5 Calsius S/I (di.. habisiMsil P

t«as

.

12) siMiN5 31 S5 -40 0 (40 50 ISO ISO SOOL b -40 -50 50 50 O I 31 31 S5 -40 0 (40 50 ISO ISO SOOL b -40 -50 50 50 O I 31 _IlI_l.Ii U.N _IlI_l.Ii U.N

ì.s.

bi

ì.s.

bi CM CISliMSISI. 0,4 CM CISliMSISI. 0,4 isiMs bi isiMs bi M s4«s 3.3 M s4«s 3.3 I. U I. U S 544.45 S 544.45

Ii

VS$

Ii

VS$ h. hil«wv, h. hil«wv, 0.5 5.I.11 0.5 5.I.11

(6)

TABLE OF CONTENTS

Page

EXECUTIVE SUMMARY _ix

LIST OF SYMBOLS _x

INTRODUCTION AND BACKGROUND 1

1.1 The Fatigue Structural Weldments 1

1.2 The Fatigue Design of Weldxnents 1

1.3 Factors Influencing the Fatigue Life of Weidments 2

1.4 Purpose of the Current Study 3

1.5 The Munse Fatigue Design Procedure (MFDP) 4

1.6 References 6

Table 7

Figures ₈

COMPARISON OF THE AVAILABLE FATIGUE LIFE

PREDICTION METHODS (TASK 1) ₁₅

2.1 Models Based on S-N Diagrams ₁₅

2.2 Methods Based upon Fracture Mechanics 18

2.3 Comparisons of Predictions with Test Results ₂₀

2.4 References ₂₁

Tables ₂₃

Figures ₂₅

FATIGUE TESTING OF SELECTED SHIP STRUCTURAL DETAILS

UNDER A VARIABLE SHIP BLOCK LOAD HISTORY (TASKS 2-4) 35

3.1 Determination of the Variable Block History 35

3.2 Development of a "Random" Ship Load History 37

3.3 Choice of Detail No. 20 and Specimen Design ₃₈

3.4 Materials and Specimen Fabrication 38

(7)

3.6 Test Results and Discussion 42

3.7 Task 2 - Long Life Variable Load History 42

3.8 Task 3 - Mean Stress Effects 42

3.9 Task 4 - Thickness Effects 44

3.10 References 47

Tables 48

Figures 64

FATIGUE LIFE PREDICTION (TASK 6) 86

4.1 Predictions of the Test Results Using the MFDP 86

4.2 Predictions of the Test Results Using the I-P Model 87

4.3 Modeling the Fatigue Resistance of Weldents 88

Tables 93

Figures 99

FATIGUE TESTING 0F SHIP STRUCTURAL DETAILS

UNDER CONSTANT AMPLITUDE LOADING (TASK 7) 112

5.1 Materials and Welding Process 112

5.2 Specimen Preparation, Testing Conditions and Test Results . 112

5.2.1 Detail No. 34 - A Fillet Welded Lap Joint 112

5.2.2 Detail No. 39-A - A Fillet Welded I-Beam

With a Center Plate Intersecting the Web

and One Flange 113

5.2.3 Detail No. 43-A - A Partial-Penetration Butt Weld 114

5.2.4 Detail No. 44 - Tubular Cantilever Beam 114

5.2.5 Detail No. 47 - A Fillet Welded Tubular Penetration 115

Tables 116

Figures 122

SUMMARY AND CONCLUSIONS 151

(8)

Page 6.2 _{The Use of Linear Cumulative Damage} _{(Task 2)}

152 6.3 _{The Effects of Mean Stress (Task} ₃₎

152

6.4 _{Size Effect (Task 4)}

153 6.5 Use of the I-P Model _{as a Stochastic Model (Task 5)}

153 6.6 _{Baseline Data for Ship Details (Task}

7) ₁₅₄

6.7 Conclusions

154

7. _{SUGGESTIONS FOR FUTURE STUDY}

156

APPENDIX A _{ESTIMATING THE FATIGUE LIFE OF WELDMENTS} USING THE IP MODEL

157 A-1 Introduction

157

A-2 _{Estimating the Fatigue Crack}

Initiation Life (N1) 158

A-2.1 _{Defining the Stress History (Task} ₁₎

158

A-2.2 _{Determining the Effects of}

Geometry (Task 2) . . 159

A-2.3 _{Estimating the Residual Stresses} _{(Task 3)}

160

A-2.4 _{Material Properties (Tasks 4-6)}

161

A-2.5 _{Estimating the Fatigue Notch}

Factor (Task 7) ₁₆₃

A-3 _{The Set-up Cycle (Task 8)}

164 A-4 _{The Damage Analysis (Task 10)}

166

A-4.1 _{Predicting the Fatigue Behavior} Under Constant Amplitude Loading With No Notch-Root Yielding or

Mean-Stress Relaxation

166

A-4.2 _{Predicting the Fatigue Behavior} Under Constant Amplitude Loading With Notch-Root Yielding and No Mean-Stress Relaxation

170

A-4.3 _{Predicting the Fatigue Crack}

Initiation Life

Under Constant Amplitude Loading with

Notch-Root Yielding and Mean-Stress Relaxation . . . . 170

A-4.4 _{Predicting the Fatigue Crack}

Initiation Life

Under Variable Load Histories Without Mean Stress Relaxation

(9)

A-5 Estimating the Fatigue Life Devoted to

Crack Propagation (Ne) 172

Figures 176

APPENDIX B DERIVATION OF THE MEAN STRESS AND THICKNESS

CORRECTIONS TO THE MIJNSE FATIGUE DESIGN PROCEDURE 196

Figure 199

APPENDIX C SCHEMATIC DESCRIPTION OF SAE BRACKET AND

TRANSMISSION HISTORIES 200

(10)

EXECUTIVE SUMMARY

This study is a continuation of a research effort at the University of Illinois at Urbana-Champaign (UIUC) to characterize the fatigue behavior of fabricated ship details. The current study evaluated the Munse Fatigue

Design Procedure and performed further tests on ship details.

The available analytical models for predicting the fatigue behavior of weidments under variable amplitude load histories were compared using test results for weidments subjected to the SAE bracket and transmission variable load amplitude histories. Models based on detail S-N diagrams such as the Munse Fatigue Design Procedure (MFDP) were found to perform well except when the history had a significant average mean stress. Models based on fatigue crack propagation alone were generally conservative, while a model based upon estimates of both fatigue crack initiation and propagation (the I-P Model) performed the best.

An extensive series of fatigue tests was carried out on welded

struc-tural details commonly encountered in ship construction using a variable load history which simulated the service history of a ship. The results from this study showed that linear cumulative damage concepts predicted the test results, but the importance of small stress range events was not

studied because events smaller than 68 MPa (10 ksi) stress range were deleted from the developed ship history to reduce the time required for

testing. _{An appreciable effect of mean stress was observed, but the results}

did not verify the existence of a specimen-size effect.

Both the Munse Fatigue Design Procedure (MFDP) and the I-P Model were

used to predict the test results. The MFDP predicted the mean fatigue life

reasonably well. Improved life predictions were obtained when the effect of

mean stress was included in the MFDP. _{Mean stress and detail size} correc-tions were suggested for the Munse Fatigue Design Procedure.

Generally good results were obtained using the I-P Model, but the

predictions for the smallest size weldments were very unconservative. The I-P model was used to develop a stochastic model for weldment fatigue

behavior based on the observed random variations in specimen geometry and induced secondary stresses resulting from distortions produced by welding. Design aids based on the I-P model are presented.

Baseline constant-amplitude S-N diagrams were developed for five

complex ship details not commonly studied in the past.

(11)

a, a., af crack length, initial crack length, final crack length

b fatigue strength exponent

C _{fatigue crack growth coefficient; also, S-N curve coefficient}

c _{fatigue ductility exponent; also, original half length of}

incomplete joint penetration and length of major axis of elliptical crack

D., Dblk

fatigue damage per cycle and per block, respectively

E Young's Modulus

F function of residual stress distribution

IJP incomplete joint penetration

K' cyclic strength coefficient

K , K fatigue notch factor and maximum fatigue notch factor

f fmax

Kt elastic stress concentration factor

Krms, KlS max min K r M M s' t' LIST OF SYMBOLS

maximum and minimum root mean square stress intensity factor

stress intensity factor due to residual stress

AK stress intensity factor range

AK root mean square stress intensity factor range

rms

k the Weibull scale or shape parameter

Ll leg length of weld perpendicular to the IJP

L2 S-N curve slope parallel to the IJP

magnification factor for free surface, width and stress

gradient

m reciprocal slope of the S-N diagram

Nf cycles to failure

2N - reversals to failure fi

N., Nfi cycles to failure at ith amplitude

(12)

n fatigue crack growth exponent; also, thickness effect

exponent

n' cyclic strain hardening exponent

ni cycles at ith amplitude

R stress ratio

RF reliability factor

r notch root radius

S, Sa remote stress range, amplitude of remote stress

remote axial stress amplitude

remote bending stress amplitude

5rms 5rms

stress and root mean square of maximum and minimum stress max min

total remote stress amplitude

Sm applied average axial mean stress

gripping bending stress

S ultimate tensile stress

u

t plate thickness; also, time

w specimen width

x ratio of remote bending stress amplitude to total remote

stress amplitude

X ratio of K to KA

max max

Y geometry factor on stress intensity factor

y(t) stress (strain) spectral ordinate for random time load history

z ratio of bending stress to axial stress

a, ß, A geometry coefficients for elastic stress concentration factor

Ef possible error in fatigue model

Sf coefficient of variation in fatigue life

LSD maximum allowable design stress range expected once during the entire life of a structure

(13)

maximum allowable design stress range expected once during

m

the entire life of a structure with allowance for baseline data and applied history mean stress

average constant amplitude fatigue strength at the desired

design life

SN R( average constant amplitude fatigue strength at the desired

design life from baseline testing conditions under stress ratio (R) = R

SN

tSN(l) average constant amplitude fatigue strength at the desired

design life from R = -1 baseline testing conditions

local strain and local strain range

fatigue ductility coefficient

e strain normal to the crack n

spectral ordinates output from the FFT analyzer (in volts)

random phase angle sampled from a uniform distribution, 02ir

a, M, a

local stress, local stress range, local stress amplitude

a , a mean stress and residual stress

r

local maximum principal stress amplitude

¿

effective residual stress amplitude fatigue strength coefficient

flank angle of welds

random load factor

uncertainty in the mean intercept of the S-N repression line

uncertainty in the fatigue data life

total uncertainty

(14)

1. INTRODUCTION AND BACKGROUND

1.1 The Fatigue Structural Weldments

Ships, like most other welded steel structures which are subjected to

fluctuating loads, are prone to metallic fatigue. While fatigue can occur

in any metal component, weldments are of particular concern because of their wide use, because they provide the stress concentrators and, because they

are, therefore, likely sites for fatigue to occur. It is for these reasons

that the fatigue of weldutents has been so exhaustively studied. However,

despite 100 years of research and thousands of studies of weldment fatigue,

there seems to be only slow progress in putting this problem to rest. This slow progress is probably due to the following:

There is a nearly infinite variety of welded joints.

Weldinents of the same j oint type are usually not exactly alike.

The behavior of even simple weldments can be exceedingly complex. The stresses in a weldment are usually imprecisely known.

The variety and complexity of the more common structural weldments are

evi-dent in Fig l-1 which shows the structural details covered in the AISC

fa-tigue provisions [l-l].

1.2 The Fatigue Design of Weldments

There are three main approaches to the fatigue design of weldments:

S-N diagrams Weldinents may be designed using the S-N curves for the

particular detail. The behavior of weldznents under constant amplitude

load-ing has been reported in the literature for hundreds of different _{j oint}

geometries. _{Attempts to collect the available information and develop} _a

weidment fatigue data base have been undertaken at _{the University of}

Illinois by Munse [l-2] and by The Welding Institute [1-3]. A typical

collection of weldznent fatigue data from the University of Illinois _Data Bank is shown in Fig. l-2 in which it is evident that the fatigue _resistance of low stress concentration fatigue-efficient weidments _{is less than plain}

plate and is characterized by a great deal of scatter. Munse [l-4] proposed a fatigue design procedure which uses the "baseline" S-N diagram information

(15)

both the desired level of reliability and the variable nature of the applied

loads (Fig. l-4). A short description of the Munse Fatigue Design Procedure

is given in Section 1-5.

Fracture Mechanics: Because fatigue is a process which begins at

stress concentrations (notches), several analytical methods of weldment

fatigue design have recently been developed which are based on mechanics

analyses of fatigue crack initiation and fatigue crack growth at the

critical locations in the structure. Such design methods or analyses

involve sophisticated, complex models (see Fig. l-4). Models based on both

fatigue crack initiation and growth have been proposed by Lawrence et al.

[l-5]: see Appendix A. Models based on fatigue crack growth alone have been

suggested by Maddox [l-6] and Shilling, et al. [l-71.

Structural Tests: A third alternative for the fatigue design of

struc-tures is to base the design on full-scale tests or observations of service

history. While such observations are closest to reality, full-scale tests

are usually prohibitively expensive and time consuming. Moreover, it is sometimes difficult to apply results from one structure to another. In the

case of ships, such tests may require a 20 year study.

1.3 Factors Influencing the Fatigue Life of Weidments

There are four attributes of weldments which, together with the magni--tude of the fluctuating stresses applied, determine the slope and intercept

of their S-N diagram: the ratio of the applied or self-induced axial and bending stresses; the severity of the discontinuity or notch which is an

inherent property of the geometry of the joint; the notch-root residual

stresses which result from fabrication and subsequent use of the welclment,

and the mechanical properties of the material in which fatigue crack

initia-tion and propagainitia-tion take place. 0f these four, the mechanical properties

are probably the least influential.

In most engineering design situations involving as-welded weldments of

a given material, the permissible design stresses are governed by: the

joint geometry, the desired level of reliability, the variable nature of_-the

(16)

indication of the sensitivity of the fatigue design stress to these design

variables. The design stress varies greatly with detail geometry, desired

level of reliability and the nature of the variable load. Mean stress has

only a modest influence.

I 1.4 Purpose of the Current Study

This report summarizes a research program sponsored by the U.S. Coast

Guard at the University of Illinois at Champaign-Urbana on the "Fatigue

Characterization of Fabricated Ship Details, Phase II" (contract DTCG

23-84-C-20018). This program is a continuation of one begun at the University of

Illinois under the direction of Professor W. H. Munse [l-4]. The second

j phase had as its principal objectives:

* To evaluate the Munse Fatigue Design Procedure developed and

dis-cussed under Phase I of the project;

* To carry out laboratory fatigue tests of fabricated ship details; * And to perform further tests on ship details.

The tasks of this study are summarized in Table 1-l.

Seven tasks were originally proposed, and they may be broken into four

categories: The first category, Task 1 was a comparison of the Munse

Fatigue Design Procedure (MFDP) predictions with the predictions resulting

from other methods of estimating the fatigue life of weldments and an

assessment of the accuracy of the Munse Fatigue Design Procedure in general.

The results of this comparison are summarized in Section 2.

In the second category, Tasks 2-4 involved long-life testing, mean

stress effects, and size effects. Each of these three tasks address a

sepa-rate issue of concern affecting our ability to predict the fatigue life of

weidments. For example, there is concern whether linear cumulative damage

is accurate in the long-life regime. Also, mean stress effects are not

generally dealt with, and there is concern that neglecting mean stress

introduces a considerable inaccuracy in the fatigue life prediction methods. Lastly, one generally ignores the influence of the absolute size of

weld-ments, and there is increasing evidence that there is an effect of size on

the fatigue life of weldinents. These phenomena were studied experimentally,

(17)

The third category was the application of the I-P Model for total

fa-tigue life prediction to the ship details considered in this program. The

I-P model was proposed as a basis for fatigue rating of ship details, but

this task (Task 5) was deleted at the outset of the program. The I-P model

in its current state of development is summarized in Appendix A. Section 4

compares the predictions made using the Munse Fatigue Design Procedure and

the I-P Model with the experimental test results (Task 6).

The fourth category (Task 7) was a program of fatigue testing of

se-lected ship details for which inadequate fatigue test data currently exists. The results of constant amplitude testing of the selected ship details is

summarized in Section 5.

1.5 The Munse Fatigue Design Procedure (MFDP)

The Munse Fatigue Design Procedure MFDP [l-41 is an effective method of design against structural fatigue and deals with the complex geometries, the

variable load histories, and the variability in these and other factors

encountered in the fatigue design of weidments.

Figure 1-2 shows the output from the University of Illinois Fatigue

Data Bank for a mild steel double-V butt weld. The Munse method fits such

data with the basic S-N relationship shown in Fig. l-3. When stress

histories other than constant-amplitude are used, different S-N diagrams

result if the test results are plotted against the maximum stress: see Fig. l-6. The Munse method accounts for this effect by introducing a term ¿

which when multiplied by the constant amplitude fatigue strength at a given

life will predict the fatigue strength for the variable load history at the

same number of cycles: see Fig. l-7.

Similarly, the natural scatter in fatigue data shown in Fig. l-8

to-gether with the uncertainties in fabrication and stress analysis are dealt

with by the MFDP through the concept of total uncertainty.

Û2 Û2 + m2 Û2

n f s c

where, Û the total uncertainty in fatigue life.

(18)

-

íCf + in which Cf is the coefficient of variation in the

fatigue life data about the S-N regression lines; and is the

error in the fatigue model (the S-N equation, including such

effects as mean stress), and the imperfections in the use of

the linear damage rule (Miner) and the Weibull distribution

approximations.

û = the uncertainty in the mean intercept of the S-N regression

lines, and includes in particular the effects of workmanship

and fabrication. A model for this uncertainty is suggested in

Section 4.3.

= measure of total uncertainty in mean stress range, including

the effects of impact and error of stress analysis and stress

determination.

Of the above mentioned sources of uncertainty, those which are best

es-timated are probably the smallest (ûf). Those which are the largest are probably the least easy to estimate (û). In modern fatigue analysis, it is

commonly believed that the greatest uncertainty is an exact knowledge of the loads to which a structure or vehicle will be subjected in service. Often the service history bears little resemblance to that which the designer

con-templates. This difficulty with application of the Munse method as with all

other design methods will require extensive field observations and

measure-ments.

Having estimated the total uncertainty in fatigue life Q, the reliabil-ity factor Rf is estimated after assuming an appropriate distribution to

characterize the load history and after specifying a desired level of

reli-ability.

The MFDP estimates the maximum allowable design stress range _{LSD from} the weldment S-N diagram by determining the average fatigue strength at the

desired design life tSN and multiplying this value by the random load

correction factor ¿ and the reliability factor (RF):

(19)

The Munse method takes all uncertainties into account and provides a

rational framework for designing structural details to a desired level of

reliability: see Fig. l-9.

1.6 References

l-l. AISC. "Specification for the Design, Fabrication and Erection of

Structural Steel for Buildings," American Institute of Steel

Construction, Nov. 1, 1978.

Radziminski, J.B., Srinivasan, R., Moore, D., Thrasher, C. and Munse, W.H. "Fatigue Data Bank and Data Analysis Investigation," Structural Research Series No. 405, Civil Engineering Studies, University of Illinois at Urbana-Champaign, June, 1973.

The Welding Institute, Proceedings of the Conference on Fatigue of Welded Structures," July 6-9, 1970, The Welding Institute, Cambridge,

England, 1971.

Munse, W.H., Wilbur, T.W., Tellalian, M.L., Nicoll, K. and Wilson, K., "Fatigue Characterization of Fabricated Ship Details for Design,"

Ship Structure Committee, SSC-318, 1983.

Ho, N.-J. and Lawrence, F.V., Jr., "The Fatigue of Weidments

Subjected to Complex Loadings," FCP Report No. 45, College of

Engineering, University of Illinois at Urbana-Champaign, Jan. 1983.

Maddox, S.J., "A Fracture Mechanics Approach to Service Load Fatigue in Welded Structures," Welding Research International, Vol. 4, No. 2,

1974.

Schilling, C.C., Klippstein, K.H., Barsom, J.M. and Blake, C.T., "Fa-tigue of Welded Steel Bridge Members under Variable Amplitude

(20)

Table 1.1

Program Summary

line fatigue resistance.

Task _Description

1. Comparison of MFC Prediction _{Compare prediction of Munse} Criterion with other predictive

methods.

2. Long Life Testing Perform long life variable load history fatigue tests on structural

details.

3. Mean Stress Effects _{Check the influence of average mean} stress on fatigue resistance under

variable load history.

4. Size Effect _{Check the influence of plate}

thickness and weld size on fatigue

resistance.

5. Fatigue Rating Deleted.

6. I-P Model Application _{Predict long life of ship structure}

through I-P Model application.

7. Fatigue Testing of Ship _{Selected structural details will be} Structural Details _{fatigue tested to determine base}

(21)

3

5Iigory

C

5 7 Full Penetrallon Il 12 3 4 18 '9 20

Fig. l-1 Structural details provided in AISC fatigue provision [l-l].

22 23 25

(22)

AW

PP r-e i. s.

---''l!.

:.

St

e UI.ullI5IpiiS1ef'iC,S -u.e!& :.

-leI. r

--un. e er

... .

_{ÒDQ seO D} D

-Mild Steel R

O AW

Butt Welds, As Welded

= Plain Plate I t

tI''

n t n n

tilt

Lower Tolerance Limit-99%

-Surv%val

-- 50% Confidence Level

-95% Confidence Level

:

Do---.

I I

iiititl

ti

I

111111!

tI

I

1111111

t I I

111111

Cycles To Failure, In Thousands

Fig. l-2

Stress range versus cycle to failure for mild steel butt welds subjected to zero to tension loading. The fatigue resistance of as-welded butt weld is generally less than the fatigue resistance of plain plate which is also indicated in

this figure [l-2]. 200 I00 80 60 u, t', ti, 40 L. _U) 8 C 4 2 2

4 68CC

2 4

681000.

2 2 4

6810

4 6 810,000

(23)

Log S

Log ii

Fig. 1-3

Basic S-N relationship for fatigue [1- 4].

Log

= Log Cm Log S

n

(24)

Determine Load Histogram

tC)etermine Structural Resoonse

Fatigue Data Bank

LS

Cycles

Determine Allowable Stress

(RF) ()

'Determine Load Histogram

t9etermine Structural Resoonse

Determine Notch-root

Stresses and Strains

Strain

(

Calculate Initiation and

I

Proragation Lives

N

I af -n

Di=l

Np-l/C

J51AK da

Fig. l-4 Fatigue design method. _{Fatigue design method based on detail}

S-N diagrams (left) such as the Munse _{approach compute the}

design stress SD based on corrections to constant amplitude fatigue resistance for the effects of variable load history

( ) and the desired reliability (RF). _{Fracture mechanics based}

design methods (right) deal with the local strain events at

the critical locations and provide _{estimates of the fatigue}

crack initiation life (N1), fatigue crack propagation life (Np) or the total fatigue life (N1 + Nr).

(25)

due to detail geometry 50% N - 10

N-106

J log cycles (N)

s-

50%

s-

50%

s-

50%

Sensitivity of S QIO level of reliability (Rel)

50% reliability - 1.00 550% 90% reliability - 0.70 95% relIability - 0.60

°'

99% reliabilIty - 0.45 S

90%

SOt

95% 99%

Sensitivity of S 0to variable load history factor:

(VLH) Welbull (k-1) S - 9.105 SO' AS R--1 Beta (q-7, p-3) S - 1.385 D SOa S

- (1 0.25R)S

50% SOt 9.9

Fig. l-5 A general indication of the sensitivity of the fatigue design stress to the design variables of

the joint geometry, the variable nature of the applied load, the desired level of reliability and the applied mean stress. Beta (q7. p7) _S - 1.895 AS D SO% Variation in S

due to mean stress

50%

I

69.4 ksl. 31.1 ksl. (100%) (100%) 20.3 ksl. 5.5 kaI. (29%) (2 1%) 21.5 kaI. 5.8 ksl. (31%) (19%) o s min. constant amplitude S - 1.005 D 50% AS Beta (q-3, p-7) S - 2.835 AS D 50%

(26)

t.

E E D' o

ij4c

laco

i':

\ I\

I\

I

/ 'I

's 's s' s-Log S 's 's

's

''s

%.s_

'

5*5

n

PI

¡P

Welded Soecimen St37/St52 I I I IO

io

lO No. of Cyclez

Fig. l-6 Fatigue resistance of a weldment subjected to variable

loadings [l-2].

n

Log n

Fig. l-7 Relationship between maximum stress range of variable (random) loading and equivalent constant-cycle stress range [l-4J. Shape of the Amplitud e

Distribution

25 20 5 I h Li. ll 108

io

(27)

Log SR

s

Constant Cycle Fatigue

- C = stm Constant Cycle. E

TL

S0 S x RF S x

(I)rn

n

Useful Mean Life

Life At For Design

L (n)

Stress

Log n

Fig. l-8 Distribution of fatigue life at a given stress level {l-4].

N

Fig. l-9 Application of reliability factor to mean fatigue resistance

(28)

2. COMPARISON OF THE AVAILABLE FATIGUE LIFE PREDICTION METHODS (TASK 1)

The effect of variable loadings on the fatigue performance of welds is

generally accounted for by using cumulative damage rules. These rules

at-tempt to relate fatigue behavior under a variable loading history to the

behavior under constant amplitude loading. The Palmgren-Miner linear

cumulative damage rule (or commonly, "Miner's rule") is widely used in many

current standards and design codes. Several models for predicting weldxnent

fatigue life have been proposed based on the S-N curve for weld details and

Miner's rule.

There are essentially two types of prediction models reported, and these are summarized in Table 2.1. The first type is based on the S-N

diagrams for the actual weld details, and the Munse Fatigue Design Procedure

is in this category. The second type is based on the fracture mechanics and

the fatigue properties of laboratory specimens, and the I-P model is in this

=1

(2-l)

where ni is the number of cycles applied at stress range ASj in the variable loading history and N is the constant amplitude fatigue life corresponding

to AS1. While Miner's rule usually gives slightly conservative life

predic-tions, it has been found to give unconservative life predictions for certain

types of variable loading history [2-l]. Two better methods of damage

accumulation have been proposed to predict the fatigue strength of weldments.

The first method uses the Miner's rule but modifies the fatigue limit

of the constant amplitude S-N curve for the welded detail. Figure 2-1 shows two typical ways of modifying the S-N curve. One way is to extend the

(29)

example, Schilling and Klippstein [2-2] have employed an equivalent stress range of constant amplitude that produces the same fatigue damage at the

variable amplitude stress range history it replaces. As the negative

reci-procal slope of S-N curve is about three for structural steel and structural details, Schilling et al. suggested the use of the "root-mean-cube (RMC)

stress range" for welded bridge details subjected to variable amplitude

loading history.

The other way suggested in BS 5400 [2-3] is changing the S-N curve from

a slope of -1/rn to -l/(m-i-2) at l0 cycles.

The second method for improving damage accumulation is to introduce a

nonlinear damage rule. In the Joehnk and Zwerneman's nonlinear damage model

[2-41, the ratio of damage to stress range increases nonlinearly as the

stress range decreases. Effective stress ranges were defined for subcycles

first, then Miner's rule was employed to calculate the damage of subcycles. Two fatigue prediction models have been proposed to predict the fatigue

resistance of welds subjected to variable loading history using constant

amplitude S-N diagram and will be discussed below: one uses Miner's rule and

an extended S-N curve, the Munse Fatigue Design Procedure, and the other uses and empirical relationship based on test results, Gurney's model.

Munse's Fatigue Design Procedure

The Munse Fatigue Design Procedure was reviewed in Section 1.5 and can

be used as a prediction method if one considers the variation in the random

variables to approach zero. Three factors are considered in Munse Fatigue Design Procedure l-4]: (a) the mean fatigue resistance of the weld

details, (b) a "random load factor" () that is a function of variable

amplitude loading history and slope of the mean S-N curve, and (e) a

"reliability factor" (RF) (roughly the inverse of the safety factor) that is

a function of the slope of the mean S-N curve, level of reliability, and a

coefficient of variation here taken to be 1.

The maximum allowable fatigue stress range SD for welds subjected to variable loading history is obtained from the following equation:

(30)

where ESN is the constant amplitude stress range at fatigue life of N

cycles. For welds subjected to a constant amplitude stress-range (SN), the

mean fatigue life N is given by the relationship:

N

(ASN)m

(2-2)

where C and m are empirical constants obtained from a least-squares analysis

of S-N diagram data. Munse's procedure uses the extended straight S-N line

at the stress ratio RO as its basis (see Fig. 2-l) and neglects the effects

of mean stress, material properties, and residual stress.

After cycle counting, the variable load history is plotted in a stress range histogram. Mean stress level and sequence effects are regarded as

secondary effects. Since random loadings for weld details usually cannot be

determined exactly, Munse's procedure uses probability distribution

func-tions to represent the weld fatigue loading. Six probability distribution

functions are employed to represent different common variable loading

histories: beta, lognormal, Weibull, exponential, Rayleigh and a shifted

exponential distribution function. It is necessary to determine which

distribution or distributions provides the best fit to a given loading

history. The random load factor in Munse's procedure are for a desired life

and are tabulated in [l-4]. Table 7.5 in [l-4] gives coefficients to adjust

values of ¿ to other design lines. In this study, the values of random load factor have been derived for any arbitrary fatigue life and are shown in

Table 2-2.

The reliability factor is given by:

[PF(N)]û°8

_)l/m Rf

-1.08 (2-3)

where PF(N) is the probability of failure, ÛN is the total uncertainty for fatigue life of N cycles and r is the gamma function.

In Ref. l-3 it is suggested that this relationship can be represented

(31)

nN.

li

N

_N([ll

N

I

b C 2 i (2-4)

where Nb = the fatigue life in blocks

= the fatigue life in cycles at maximum stress range in the block

his tory

Ni = number of cycles per block equal or exceeding Pj times the maxi-mum stress range in the block history

n = total number of cycles in a block

The parameter contained within the braces is the random load factor.

2.2 Methods Based upon Fracture Mechanics

Methods based upon fracture mechanics ignore the fatigue crack

initia-tion phase and calculate the fatigue crack propagainitia-tion life only. Maddox

[l-61

used linear fracture mechanics and Miner's rule to predict the fatigue life of welds subjected to variable loading history. Miner's rule was found to be accurate for welds under loading histories without stress interaction.

Barsom

[l-61

used a single stress intensity factor parameter,

root-mean-square stress intensity factor, to define the crack growth rate under

both constant and variable amplitude loadings. The root-mean-square stress

intensity factor AK , is characteristic of the load distribution and is rms

independent of the order of the cyclic load fluctuations. Hudson [2-6]

applied the root-mean-square (RMS) method for random loading history with

variable minimum load. This simple RMS approach has been shown applicable

50% Reliability _RF l_00

90% Reliability RF - 0.70

95% Reliability _RF 0.60

99% Reliability RF = 0.45

Gurney's Model

Gurney [2-5 J performed fatigue tests on fillet welded joints using

simple variable loading history. It was found that the logarithm of number of blocks to faílure varied linearly with the ratio of the subcycle's stress

(32)

for loading history with random sequences. The root-mean-square stresses

are defined as:

and rrnS ₌ (S )211/2 max LN max _J n=l N = (Sa. )2]l/2 min N min n=l (2-5) (2-6)

where S and S . are the maximum and minimum stress for each cycle

max min

respectively, and N is the total number of cycles for the random loading

history.

The root-mean-square stress intensity factor range is calculated from

AK = KrmS

-rms max min (2-7)

Calculation of fatigue crack propagation life is through the substitution of Eq. 2-7 into the fatigue crack propagation model, Eq. A-18.

A deterministic model for estimating the total fatigue life of welds

has been developed by the authors and is presented in Appendix A. This

model is termed the initiation-propagation (I-P) _{or total life model and}

assumes that the total fatigue life of a weld (NT) is composed of a fatigue

crack initiation (N1) and a fatigue crack propagation period (Np) such that:

NT N1 + Np (2-8)

The initiation portion of life may be estimated using the fatigue data from strain-controlled fatigue tests on smooth specimens. The initiation

life so estimated includes a portion of life which is devoted to the

development and growth of very small cracks. The fatigue crack propagation

portion of life may be estimated using fatigue crack propagation data and an

arbitrarily assumed initiated crack length (ai) of O.01-in. in the instances in which the initial crack length is not obvious. A second alternative is

to assume that a is equal to ah the threshold crack length. In most

(33)

within a factor of 2 [1-5]. Naturally, for welds containing crack-life

defects, N1 may be very short. However, for other internal defects having

low values of Kt such as slag or porosity, N1 may be appreciable; and

neglecting N1 may be overly conservative. This is particularly the case for welds containing no discontinuities other than the weld toe. In this case

and particularly for the long life region, it is believed that the fatigue

crack initiation portion life (as defined) is very important. A detailed

discussion of the I-P model is given in Appendix A.

2.3 Comparisons of Predictions with Test Results

Table 2.1 sununarizes the prediction models discussed above. Several of

these models were used to predict the "mean fatigue lives" of welds tested

in this and other studies [2-7]. Figures 2-2 to 2-10 compare the predic-tions made by the Munse Fatigue Design Procedure, Miner's rule, Gurney's model, the RMS method, and the I-P model with actual test data for several

histories. The Munse Fatigue Design Procedure (MFDP) and the Miner's Rule

predictions in these figures differ only in that the MFDP uses a continuous

probability distribution function to model the load history while the

Miner's rule sums the actual history. The "Rainflow" counting method was

used in these comparisons. In these comparisons, the maximum stress in the

load history (5Arn or S ) is plotted against the predicted life. The

min max

effects of bending stresses were taken into account.

The Munse Fatigue Design Procedure (MFDP) provided good mean fatigue

life predictions for welds subjected to the SAE bracket history (See

Appendix C) as shown in Figs. 2-2, 2-3, and 2-5. For welds tested under the SAE transmission history (See Appendix C), unconservative predictions were

made by the MFDP (Fig. 2-4). This discrepancy might be due to means stress effects because the transmission history has a tensile mean stress while the

bracket history has only a small average mean stress. The root-mean-square method (fatigue crack propagation life only) gave conservative predictions

for all cases. It is interesting to note that the predictions made based on

S-N curves without cutoff and Miner's rule are similar to the predictions of

the MFDP. Predictions resulting from the Total Fatigue Life (I-P) model seem to agree well with the test results. Table 2-3 is a statistical

(34)

sum-mary of the departures of predicted lives from the test data as in Fig. 2-6

to Fig. 2-10.

While the agreement between the prediction methods discussed above and the two variable load histories employed in the comparison _{are quite good,} there are histories for which all predictions methods based _{on linear} cumu-lative damage fall short even when the very conservative assumption of an

extended S-N diagram is used [2-8]. These histories are typically very long histories in which most of the damaging cycles are near the constant

amplitude S-N diagram endurance limit. Neither the SAE bracket or

transmis-sion histories nor the edited history discussed in the next section fall into this category; consequently, this serious problem in fatigue life

prediction is not addressed by the comparison of this section nor the

experimental study of the next section.

2.4 References

2-l. _{Fash, J.W., "Fatigue Life Prediction for Long Load Histories," Digital}

Techniques in Fatigue, S.E.E. mt. Conf., City University of _London,

England, March 28-30, 1983, pp. 243-255.

2-2. _{Schilling, C.C. and K.H. Klippstein, "Fatigue of Steel Beams} _by

Simu-lated Bridge Traffic," Journal of Structural Division, Proceedings of

ASCE, Vol. 103, No. ST8, August, 1977.

2-3. _{BS5400: Part lO: 1982, "Steel Concrete and Composite Bridges,} _{Code of} Practice of Fatigue."

2-4. Zwerneman, F.J., "Influence of the Stress Level of Minor Cycles _on

Fatigue Life of Steel Weldments," Dept. of Civil _{Engineering, The}

University of Texas at Austin, Master Thesis, May 1983.

2-5. _{Gurney, T.R., "Some Fatigue Tests on Fillet Welded Joints under Simple}

Variable Amplitude Loading," The Welding Institute, May 1981.

2-6. Hudson, C.M., _{"A Root-Mean-Square Approach for Predicting Fatigue}

Crack Growth under Random Loading," ASTM STP 748, 1981, _{pp. 41-52.}

2-7. Yung, J. -Y. and Lawrence, "A Comparison of Methods for Predicting

Weldment Fatigue Life under Variable Load Histories," FCP _{Report No.} 117, University of Illinois at Urbana-Champaign, Feb., _1975,

2-8. Gurney, T.R., _{"Fatigue Test on Fillet Welded Joints to Assess the} Validity of Miner's Cumulative Damage Rule," Proc. Roy. _{Soc., A386,}

(35)

2-9. Miner, M.A., "Cuxnalative Damage in Fatigue," Journal of Applied

(36)

Basis Proposed by

S-N curve Miner [2-9]

Summary of Fatigue Life Prediction Models For Weidments Subjected to Variable Loadings

Zwerneman [2-4] Joehnk Gurney [2-5] Munse [l-4] fracture Barsom [l-7] mechanics Lawrence [l-5] Ho Table 2.1 Model (ni/Ni) = 1

ni : no. of cycles applied at

no. of cycles to failure at linear damage accumulation

ASeff ASi(ASmax/ASi)a

ASeff : effective stress range at ASmax

AS1 : stress range of subcyles

ASmax : maximum stress range

a : varies with loading history

nonlinear cumulative damage

n

p-N N [ fi (N . /N .)

'J

b c

2 ei-i ei

Nb : no. of blocks to failure

N : no. of cycles to failure at ASmax

Nei: no. of cycles per block equal to or exceeding Pi times the maximum stress in one block

SD - SN * * RF

SD : allowable maximum stress range

SN : maximum stress range in life N

probabilistic random load factor reliability factor

rms = [( A}()2/n]1/2

rms : root mean square stress intensity

factor range

fatigue crack propagation life only

NT N1 + N

NT : total Eatigue life

N1 : fatigue crack initiation life

(37)

Distribution Function Random Load Factor, ¿ beta ([r(q)r(m+qr)]/[1'(m+q)l'(q+r)]) Weibull (inN)hh'k[r(l+m/k)] -1/rn exponential (inN) [r(l+m)]*1/rn Rayleigh

(jp)l/2[r(l+m/2)]/m

lognormal shifted exponential Û : Coefficient of Variation of F Fp P Table 2.2

Random Load Factors for Distribution Functions [1-4]

i m [ m!/(mn)!(2nNYn(l

-a)a

nm-n -1/rn n=0 a ₌ _{a/[a-4-z(inNb))} Table 2.3

Statistical Summary of the Departures of Predicted Lives from Fatigue Test Data

1/rn No. of Cases 29 29 29 13 29 1.061 1.015 0.894 0.906 1.016 cl 0.124 0.093 0.081 0.052 0.067 Fp log10 (N _{di ti} Mean value of F; F

log10

unity nf F

value

Test'

represents the perfect agreement between the prediction and fatigue

data.

Munse' s Miner' s Curney' s RMS Method I-P Model

(38)

S-N Curve (In Air)

m

f

With Fatigue Limit

rBS 5400

J

m+2

Extended Line

-Log Nf, cycles

Fig. 2-1 Modification of S-N diagram.

(39)

Io'

I

Il

GM MS4361 Steel

Fillet Welds

SAE Bracket History

IJP Failure

-: Test Results

:1-P Model

-- :

Muns&s Beta (S-N)

--- :

Miner's Rule (S-N)

RMS Method

I I i i

till

'a'

IO'

102

NT,

blocks

Fig. 2-2

Fatigue test results and predictions for GM MS4361 Fillet welded cruciform joints subjected to SAE Bracket history.

S1

is the minimum stress in the load history.

The Incomplete Joint penetration (IJP) is indicated by 2C.

I I I I I

titi

I I I I

titi

I I t I I

IIi

(40)

102 lo'

CT 1E650-B Steel

Fillet Welds

-SAE Bracket History

.Ijp Failure

A

: Test Results

I-P Model

-- :

Munse's Beta (S-N)

--.- : Miner s Rule (S-N)

RMS Method

i I I I i

Iii

I I ¡ I I

IIIj

-A

I I I I I

lIti

4

IO'

102

IO IO4

NT,

blocks

Fig. 2-3

Fatigue test results and predictions for CT 1E650-B

fillet welded cruciform joints

subjected to SAE Bracket history.

Smtn is the minimum stress in the load history.

The Incomplete Joint penetration

(41)

ASTM A36 Steel

Butt Welds

SAE Transmission History

Toe Failures

A

:

Test Results

I-P Model

--: Munse's Rayleigh (S-N)

Miner's Rule (S-N)

--: RMS Method

__P

I I I

111111

I I I

IIIlIJ

I 1

111111

D

IO'

X D

<E

(J)

lo'

102 IO3

NT

blocks

Fig. 2-4

Fatigue test results and predictions for ASTM

A 36

butt welds subjected to SAE

transmission

history

SA

is the maximum stress in the

load

history.

min

(42)

102

ASTM A588-A Steel

Butt Welds

SAE Bracket History

Toe Failure

: Test Results

I-P Model

--: Munse's Beta (S-N)

-.-: Miner's Rule (S-N)

N1, blocks

Fig. 2-5

Fatigue test results and predictions for ASTM A588-A butt welds subjected to SAE Bracket history.

S

is the minimum stress in the load history.

20 10û

lO'

I-

C

<E

w

I&

102

l0

(43)

102

I I I

IIIIt

i J

1111111

I I

I'IIII

:

Cruciform Joints

/

:GMMS436I

/

_/1

-

A:CTIE65O-B

/

-Butt Welds

/

o: ASTM A588-A

/

: ASTM A36

/

AAA/

/

/.

A

_/

/

0/

_/

/

s

/

Munses Fatigue Criterion

i

iii iii

i i i

ii iii

i i i i

t tu

/

Actual Life, blocks

Fig. 2-6 Comparison of actual and predicted fatigue life using Munse's

Fatigue Criterion (Extended S-N Curve). The dashed lines

(44)

u,

O ir4

o

-Q

/

Cruciform Joints

e: GM MS4361

A

CT lE650-B

Butt Welds

o: ASTM A588-A

: ASTM A36

/

_/

/

=

.,

_/

0/

A

/

Miner1s Rule

-,/'

_{(Based on Straight}

2'

S-N Curves)

ic

i

j titil

i i i

ill iii

I I I

II

11

102

_IO

IO4

_IO5

Actual Life, blocks

Fig. 2-7 _{Comparison of actual and predicted fatigue life using Miner's}

Rule and the Extended S-N Curve. The dashed línes represented

(45)

102

-

I I I

111111

:

Cruciform Joints

/

- : GM MS4361

=

:CTlE65O-B

/

Butt Welds

0:

ASTM A588-A

A:

ASTM A36

/

-

/

I

/

-

/

-

/

-

/

_/0

/

A

/A

°

Gurneys Model

/

¿

I I

i III

I I I I

li iii

i i t I

lilt

Actual Life, blocks

I I I i

Ili5

¡ i

III

Fig. 2-8 Comparison of actual and predicted fatigue life using Gurney's

Model. The dashed lines represented factors of two departures

(46)

a)

-J

-o

a)

.4-o

-o

a) û-I I I

I II

I I I I I I I j I I ¡ I

Cruciform Joints

/

s: GM MS4361

//

_/

£:CTIE65O-B

_/

/

_/

/

_/

/

_/

/

_/

/

_/

/

_/.

io3

/

/AA

A

//

-

/

A

//

RMS Method

//

_/

/s

/

102 i

jiltilil

i

IIIHII

102

io3

Actual Life, blocks

Fig. 2-9 _{Comparison of actual and predicted fatigue life using the}

RMS Method. _{The dashed lines represented factors of two}

(47)

102

102 ìctual Life, blocks

Fig. 2-10 Comparison of actual and predicted fatigue life using the

t-P model. The dashed lines represented factors of two

(48)

3. FATIGUE TESTING OF SELECTED SHIP STRUCTURAL DETAIL

UNDER A VARIABLE SHIP BLOCK LOAD HISTORY (TASKS 2-4)

The experimental portions of this study can be divided into two parts:

The first and major effort of this study was to test a selected structural

detail under a variable load history which simulated a ship history. This

part encompassed Long Life Variable Load Testing (Task 2), Mean Stress

Ef-fects (Task 3), and Thickness Effects (Task 4). The results of the first part of the experimental program are discussed in this section. The second

part of the experimental program (Task 7) was the collection of baseline

constant amplitude fatigue data for selected ship structural details. The

results of this latter study are summarized in Section 5.

A major technical difficulty at the outset of this part of the

experi-mental program was obtaining a variable load history which simulated the

typical service history experienced by ships. Since no standard ship

his-tory was available, the first major task for the experiments described in

this section was to develop a reasonable variable load history block which

simulated the load history of a ship. This task was further complicated by the fact that typical ship histories have occasional large overloads which

cannot be contained in every block or repetition of a short history and by

the large number of small cycles which contribute little to the accumulation of fatigue damage but which enormously increase the time required for

test-ing and consequently determine whether or not the laboratory testtest-ing can be

completed in a reasonable time.

3.1 Determination of The Variable Block Load History.

The Weibull distribution was demonstrated to be an appropriate

proba-bility distribution for long-term histories through comparison with actual

data such as the SL-7 container ship history [3-l]: see Fig. 3-l.

Munse [l-4] used the 36,011 scratch SL-7 gauge measurements or records taken over four-hour periods shown in Fig. 3-1. Each measurement or short

history contained 1,920 cycles. The biggest "grand cycle" of each history

was termed an occurrence, and the 36,011 occurrences were assembled into the

histogram shown in Fig. 3-2 and fitted with a Weibull distribution (k = 1.2,

(49)

histogram for the entire history composed of 52,000 short histories containing 1,920 cycles each, or io8 cycles. Using the fitted Weibull

distribution, Munse estimated the maximum stress range expected during the

ship life of 108 cycles (S10-8) by assuming that the probability associated

with this stress range would be 1/108, that is, equal to that for the

largest occurrence. From this argument and the fitted Weibull distribution,

a maximum stress range of 235 MPa (34.11 ksi) was calculated as the maximum

stress in the 20 year ship life history for the location at which the stress

history was recorded.

The SL-7 history and Munse's Weibull distribution representation of it

was adopted for use in this study. The next problem was to create a typical

history, that is a sequence of stress ranges which represented the typical

ship experience (period of normal sea state interdispersed with storm

epi-sodes), which conformed to the overall Weibull distribution. Furthermore,

to permit long-life fatigue testing, the history had to be edited to remove

cycles which caused little fatigue damage but needlessly extended the

required testing time. It was decided to edit the history so that one

"block" would contain only 5,047 cycles and yet contain the most damaging

events in a typical one month (345,600 cycle) ship history (see Fig. 3-6). The first step was to decide which of the events in the SL-7 history

were the most damaging and which were the least damaging and could therefore

be omitted. The damage calculated for a given interval of stress range of

the SL-7 history depends upon three things: the method of summing damage (linear accumulative damage or Miner's rule was used); the assessment of damage caused by a given stress range (we used the I-P model [l-51 rather

than the extended S-N approach of Munse and this makes a big difference in

what can be omitted); and the degree of stress concentration by the weld defects to be studied (we assumed a maximum fatigue notch factor (K )

fmax

typical for Detail No. 20 as K = 4.9). fmax

The justification for adhering to the predictions of our I-P model is

that it has given reasonable estimates of weldment fatigue life under

varia-ble load histories in laboratory air [2-6]: see Fig. 2-10. A major

differ-ence between the I-P model and Munse's approach is the anticipated behavior

of the weidments in the long-life region. Fig. 3-3 shows the extended line

(50)

exaggerates the importance of the smaller stress ranges and leads to the

conclusion that they can not be deleted. The I-P model predicts that the S-N curve has a slope of about 1:10 in the long life region and consequently,

predicts a lesser importance for the smaller stress ranges: see Fig. 3-4.

We used the following strategy for editing the SL-7 history. _{If each} cycle had an average period of 7.5 seconds as reported, a one month ship

history would consist of 345,600 cycles [3-fl. Keeping only those stress

ranges which contributed 92.8% of the total damage (estimated using the I-P model and Miner's rule, see Figs. 3-4 and 3-5) leads _{to the elimination of}

stress range less than 68.9 MPa (10 ksi) and greater than 152 MPa (22 ksi). This decision would permit a reduction in length of the one month history

from 345,600 (total) cycles to 5,047 cycles. _{Fig. 3-6 shows the developed}

"one-month history" which starts with a period of low stress range, 75.8 MPa (11 ksi), and gradually increases to a maximum of 145 MPa (21 ksi) during the central storm period after which the amplitude decreased _{to the original}

11 ksi. At a testing frequency of 5 Hz, a block required about 17 _minutes.

Since a 5,047 cycle block represents 345,600 cycles in _{service, 290 blocks}

or 3.5 days of testing at 5 Hz or 10 to 15 days at lower testing _frequencies are equivalent to a 108 _{cycle service history or 24 years of service.}

The ship block load history shown in Fig. 3-6 was read into the memory of a function generator which controlled _{a 100 kip MTS} _{fatigue testing}

machine.

3.2 Development of a "Random" Ship Load History

At the suggestion of the advisory committee, _{an alternative "random"}

time history was generated using a method employed by _{Wirsching [3-2].} _To simulate a stress history from a given spectral density function, _the

spec-tral density must be discretized. _{This operation was accomplished by}

defin-ing n random frequency intervals, _{in the region of definition of f.} The value of _{must be random to insure that the simulated process is a} nonperiodic function, f is the midpoint of iïf1. _{The simulation is}